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4-1 CHAPTER 4 EVALUATION OF HAZARDS ASSOCIATED WITH SOIL  LIQUEFACTION AND GROUND FAILURES Soil Liquefaction Hazards Introduction Experience from past earthquakes demonstrates the vulnerability of waterfront structures and lifelines to seismically induced ground deformations. In an extensive review of the seismic  performance of ports, Werner and Hung (1982) concluded that by far the most significant source of earthquake damage to waterfront structures has been pore pressure build-up in loose to medium-dense, saturated cohesionless soils that prevail in coastal and river environments. This observation has been supported by the occurrence of liquefaction-induced damage at numerous  ports in the past decade (e.g., Seed et al. 1990; Chung, 1995; Mejia et al., 1995; Werner and Dickenson, 1996). Components of marine facilities conspicuous for poor performance due to earthquake-induced liquefaction include pile supported structures, sheet-pile retaining walls and  bulkheads, and gravity retaining walls founded on or backfilled with loose sandy soils. The observed damage patterns commonly reflect the deleterious effects of (a) poor foundation soils which may have marginal static stability and which also tend to amplify the strong ground motions at these sites, and (b) the combination of high ground water levels and the existence of very loose to medium dense, sandy backfill and foundation soils. Saturated, sandy soils are susceptible to earthquake-induced liquefaction , a state wherein excess pore water pressures generated in the soil result in a temporary reduction, or complete loss, of strength and stiffness of the soil. The liquefaction of a loose, saturated granular soil occurs when the cyclic shear stresses/strains passing through the soil deposit induce a progressive increase in the pore water  pressure in excess of hydrostatic. During an earthquake the cyclic shear waves that propagate upward from the underlying bedrock induce the tendency for the loose sand layer to decrease in volume. If undrained conditions during the seismic disturbance are assumed, an increase in pore water pressure and resulting decrease of equal magnitude in the effective confining stress is required to keep the loose sand at constant volume. The degree of excess pore water pressure generation is largely a function of the initial density of the sand layer and the magnitude and duration of seismic shaking. In loose to medium dense sands pore pressures can be generated which are equal in magnitude to the confining stress. At this state, no effective (or intergranular) stress exists between the sand grains, and a complete loss of shear strength is temporarily experienced. The potential for the development of large strain (or flow) behavior is controlled by the initial relative density of the soil. The phenomena associated with the loss of strength of the sandy soils (e.g.; loss of bearing capacity, lateral spread, increase in active lateral earth pressures against retaining walls, loss of passive soil resistance below the dredge line and/or adjacent to anchor systems, and excessive settlements and lateral soil movements) contribute to the excessive deformations of waterfront structures. The large deformations associated with the failure of waterfront retaining walls can result in damage to wharf and backland, adversely affecting the operation of marine oil terminals.

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CHAPTER 4 EVALUATION OF HAZARDS ASSOCIATED WITH SOIL

 LIQUEFACTION AND GROUND FAILURES

Soil Liquefaction Hazards

Introduction

Experience from past earthquakes demonstrates the vulnerability of waterfront structures

and lifelines to seismically induced ground deformations. In an extensive review of the seismic

 performance of ports, Werner and Hung (1982) concluded that by far the most significant source

of earthquake damage to waterfront structures has been pore pressure build-up in loose to

medium-dense, saturated cohesionless soils that prevail in coastal and river environments. This

observation has been supported by the occurrence of liquefaction-induced damage at numerous

 ports in the past decade (e.g., Seed et al. 1990; Chung, 1995; Mejia et al., 1995; Werner and 

Dickenson, 1996). Components of marine facilities conspicuous for poor performance due to

earthquake-induced liquefaction include pile supported structures, sheet-pile retaining walls and 

 bulkheads, and gravity retaining walls founded on or backfilled with loose sandy soils. The

observed damage patterns commonly reflect the deleterious effects of (a) poor foundation soils

which may have marginal static stability and which also tend to amplify the strong ground 

motions at these sites, and (b) the combination of high ground water levels and the existence of 

very loose to medium dense, sandy backfill and foundation soils. Saturated, sandy soils are

susceptible to earthquake-induced liquefaction, a state wherein excess pore water pressures

generated in the soil result in a temporary reduction, or complete loss, of strength and stiffness of 

the soil.

The liquefaction of a loose, saturated granular soil occurs when the cyclic shear 

stresses/strains passing through the soil deposit induce a progressive increase in the pore water 

 pressure in excess of hydrostatic. During an earthquake the cyclic shear waves that propagate

upward from the underlying bedrock induce the tendency for the loose sand layer to decrease in

volume. If undrained conditions during the seismic disturbance are assumed, an increase in pore

water pressure and resulting decrease of equal magnitude in the effective confining stress is

required to keep the loose sand at constant volume. The degree of excess pore water pressure

generation is largely a function of the initial density of the sand layer and the magnitude and 

duration of seismic shaking. In loose to medium dense sands pore pressures can be generated 

which are equal in magnitude to the confining stress. At this state, no effective (or intergranular)

stress exists between the sand grains, and a complete loss of shear strength is temporarily

experienced. The potential for the development of large strain (or flow) behavior is controlled bythe initial relative density of the soil. The phenomena associated with the loss of strength of the

sandy soils (e.g.; loss of bearing capacity, lateral spread, increase in active lateral earth pressures

against retaining walls, loss of passive soil resistance below the dredge line and/or adjacent to

anchor systems, and excessive settlements and lateral soil movements) contribute to the

excessive deformations of waterfront structures. The large deformations associated with the

failure of waterfront retaining walls can result in damage to wharf and backland, adversely

affecting the operation of marine oil terminals.

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One pertinent example,  Figure 4-1,  is provided by the observations made at the U.S.

 Naval Station at Treasure Island after the 1989 Loma Prieta earthquake. Inspection of the

acceleration record obtained at the Treasure Island Fire Station shows that at about 15 seconds

after the start of recording, the ground motion was subdued; this was probably caused by the

occurrence of subsurface liquefaction. Liquefaction occurred after about 4 or 5 “cycles” of 

shaking, about 5 seconds of strong motion. Sand boils were observed at numerous location and  bayward lateral spreading occurred with associated settlements. Ground cracking was visible

with individual cracks as wide as 6 inches. Overall lateral spreading of 1 foot was estimated.

Ground survey measurements indicate that settlements of 2 to 6 inches occurred variably across

the island and that some areas had as much as 10 to 12 inches of settlement. The liquefaction

related deformations resulted in damage to several structures and numerous broken underground 

utility lines, Egan (1991).

Damping = 5%

Surface

Rock

Figure 4-1: Soil Response at Treasure Island During the 1989Loma Prieta Earthquake (After Seed et al., 1990)

Evaluation of Liquefaction Susceptibility

Experience from liquefaction-induced damage to structures and lifelines during past

earthquakes shows that liquefaction hazards can be broadly classified into three general modes:

(a) global instability and lateral spreading; (b) localized liquefaction hazard; and (c) failure or 

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excessive deformation of walls and retaining structures. These liquefaction hazards may vary

dramatically in scale, and the extent of each of these potential soil failure modes must be

evaluated for projects in seismically active regions. The scope of the investigation required will

reflect the nature and complexity of the geologic site conditions, the economics of the project,

and the level of risk acceptable for the proposed structure or existing facility.

The evaluation of liquefaction hazard is generally performed in several stages that include:

(a) preliminary geological/geotechnical site evaluation; (b) quantitative evaluation of liquefaction

 potential and its potential consequences; and, if necessary (c) development of mitigation and 

foundation remediation programs. A generalized flow chart for the evaluation of liquefaction

hazards to pile supported structures is presented in Figure 4-2. This simplified chart is intended 

to illustrate the basic procedures involved in evaluating potential liquefaction hazards and 

developing mitigation programs.

Preliminary Site Investigation  A preliminary site evaluation may involve establishing the

topography, stratigraphy, and location of the ground water table at the project site. Thesegeologic site evaluations must address the following three basic questions:

a)  Are potentially liquefiable soil types present?

 b)  Are they saturated, and/or may they become saturated at some future date?

c)  Are they of sufficient thickness and/or lateral extent as to pose potential risk with respect to

major lateral spreading, foundation bearing failure or related settlements, overall site

settlements, localized lateral ground movements, or localized ground displacements due to

“ground loss”?

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NO NO

Figure 4-2: A Flow Chart For The Evaluation Of Liquefaction

Hazards To Pile Supported Structures

The general geologic information, combined with ground motion data related to the

design-level seismic events, can be used to provide a preliminary indication of the potentialliquefaction susceptibility of soils at the site. This methodology has been developed by Youd and 

his co-workers, and is described in a recent state-of-the-art paper on the mapping of earthquake-

induced liquefaction (Youd, 1991). In addition, California Department of Conservation, Division

of Mines and Geology has established guidelines for mapping areas which might be susceptible

to the occurrence of liquefaction. These zones establish where site-specific geotechnical

investigations must be conducted to assess liquefaction potential and, if required, provide the

technical basis to mitigate the liquefaction hazard. The following is taken directly from their 

criteria:

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Liquefaction Hazard Zones are areas meeting one or more of the following

criteria:

1. Areas known to have experienced liquefaction during historic earthquakes.

Field studies following past earthquakes indicate liquefaction tends to recur atmany sites during successive earthquakes

2. All areas of uncompacted fills containing liquefaction susceptible material that

are saturated, nearly saturated, or may be expected to become saturated.

3. Areas where sufficient existing geotechnical data and analyses indicate that the

soils are potentially liquefiable. The vast majority of liquefaction hazard areas are

underlain by recently deposited sand and/or silty sand. These deposits are not

randomly distributed, but occur within a narrow range of sedimentary and 

hydrologic environments. Geologic criteria for assessing these environments are

commonly used to delineate bounds of susceptibility zones evaluated from other criteria, such as geotechnical analysis (Youd, 1991) . Ground water data should be

compiled from well logs and geotechnical borings. Analysis of aerial photographs

of various vintages may delineate zones of flooding, sediment accumulation, or 

evidence of historic liquefaction. The Quaternary geology should be mapped and 

age estimates assigned based on ages reported in the literature, stratigraphic

relationships and soil profile descriptions. In many areas of Holocene and 

Pleistocene deposition, geotechnical and hydrologic data are compiled.

Geotechnical investigation reports with Standard Penetration Test (SPT) and/or 

Cone Penetration Test (CPT) and grain size distribution data can be used for 

liquefaction resistance evaluations.

4. Areas where geotechnical data are insufficient. The correlation of Seed et al.

(1985) , and the (N1)60 data can be used to assess liquefaction susceptibility. Since

geotechnical analyses are usually made using limited available data the

susceptibly zones should be delineated by use of geologic criteria. Geologic cross

sections, tied to boreholes and/or trenches, should be constructed for correlation

 purposes. The units characterized by geotechnical analyses are correlated with

surface and subsurface units and extrapolated for the mapping project.

CDMG criteria uses the minimum level of seismic excitation for liquefaction hazard zones to be

that level defined by a magnitude 7.5-weighted peak ground surface acceleration for UBC S2

soil conditions with a 10 percent probability of exceedance over a 50-year period.

In areas of limited or no geotechnical data, susceptibility zones are identified by CDMG

geologic criteria as follows:

(a) Areas containing soil deposits of late Holocene age (current river channels and 

their historic floodplains, marshes and estuaries), where the magnitude 7.5-

weighted peak acceleration that has a 10 percent probability of being exceeded in

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50 years is greater than or equal to 0.10 g and the water table is less than 40 feet

 below the ground surface; or 

(b) Areas containing soil deposits of Holocene age (less than 11,000 years), where

the magnitude 7.5-weighted peak acceleration that has a 10 percent probability of 

 being exceeded in 50 years is greater than or equal to 0.20 g and the historic highwater table is less than or equal to 30 feet below the ground surface; or 

(c) Areas containing soil deposits of latest Pleistocene age (between 11,000 years

and 15,000 years), where the magnitude 7.5-weighted peak acceleration that has a

10 percent probability of being exceeded in 50 years is greater than or equal to

0.30 g and the historic high water table is less than or equal to 20 feet below the

ground surface.

According to CDMG, the Quaternary geology may be taken from existing maps, and hydrologic

data should be compiled. Application of this criteria permits development of liquefaction hazard maps which definite regions requiring detailed investigation, allowing concentration of sampling

and testing in areas requiring most delineation.

Quantitative Evaluation of Liquefaction Resistance If the results of the preliminary site

evaluation indicate that more in-depth studies are warranted, then additional geotechnical

characterization of the soils will be necessary. Guidelines for the analysis and mitigation of 

liquefaction hazards have been presented by the CDMG (Special Publication No. 117). In the

context of a factor of safety, the occurrence of liquefaction can be thought of as the capacity of 

the soil to resist the development of excess pore pressures versus the demand imposed by the

seismic ground motions. In practice, the quantitative evaluation of liquefaction resistance is

usually based on in-situ Standard Penetration Test (SPT) data (Seed and Idriss, 1982; Seed and De Alba, 1986) and/or Cone Penetration Test (CPT) data (Robertson et al., 1992). The

liquefaction resistance of a sand is related to the penetration resistance obtained by either the SPT

(N in blows/30 cm (or foot)) or CPT (q c in kg/cm2 (or tsf)). The penetration values measured in

the field are corrected to account for confining stresses and normalized to obtain a value which

corresponds to the N-value that would be measured if the soil was under a vertical effective

stress of 1 kg/cm2 (1 tsf). Additional corrections may be required for SPT N-values depending on

the type of drive hammer and release system, length of drill stem, and other factors as outlined by

Seed and Harder (1990). The corrected N-value used in the liquefaction analysis is designated 

(N1)60.

The earthquake-induced cyclic stresses in the soil can be estimated by (a) the simplified 

evaluation procedure developed by Seed and Idriss (1982), or (b) performing a dynamic soil

response analysis. The Seed and Idriss technique yields an estimate of the ratio of the average

cyclic shear stress on a horizontal plane in the soil to the initial vertical effective stress on that

 plane. Correlations of the penetration resistance of soils and cyclic stress ratio (CSR) at sites

which did or did not liquefy during recent earthquakes have been established for level ground 

conditions. In practice, the cyclic stress ratio (CSR) required for the “triggering” of liquefaction

can be determined once the penetration resistance of the soil has been obtained by the use of 

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liquefaction boundary curves, Figure 4-3). The cyclic stress ratio developed during an earthquake

is given by the equation:

( ) 0.65a

g

r r av

v

v

v

d MSF  

τ 

σ 

σ 

σ ′ ′

≈ ⋅ ⋅ ⋅ ⋅max

where τ av  is the average cyclic shear stress, σv' is the vertical effective stress, σv  is the total

vertical stress, amax is the maximum horizontal ground surface acceleration, g is the acceleration

of gravity, r d  is a stress reduction factor which accounts for the fact that the soil column above

the soil element behaves as a deformable body, and r MSF is the magnitude scaling factor which is

used to convert the CSR required for liquefaction due to a magnitude 7.5 earthquake (the basis

for the boundary curves in Figure 4-3 to the magnitude of interest.

a) b)

Figure 4-3:Liquefaction Boundary Curves;

A) Seed Et Al., 1975, B) After Mitchell And Tseng, 1990B) 

Several practical points should be noted regarding these plots:

a)  The N-values used in the development of the relationship were obtained using an ASTM

standard sampler driven by a 64 kg (140 lb) weight falling 76 cm (30 in). In light of the

variety of soil samplers and driving mechanisms (e.g.; safety hammers, donut hammers, slip-

 jars, etc.) commonly used in practice, the engineer should realize that correlations between N-

values obtained by various methods are tenuous at best. Appropriate caution should be

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exercised when interpreting potential liquefaction behavior based on N-values from non-

standard techniques.

 b)  The boundaries between liquefaction and non-liquefaction are based on case histories for the

surface-evidence of liquefaction. It is possible that sites classified as not exhibiting

liquefaction experienced the development of significant excess pore pressures that were notmanifested at the surface. The boundaries are therefore not intended to delimit the definitive

occurrence or nonoccurrence of liquefaction, but rather an approximate indication of whether 

ground failures may be experienced.

c)  The intensity of the ground motions are accounted for in the formulation of the CSR. The

number of load cycles is also an important parameter. To represent the effects of the number 

of cycles, a magnitude scaling factor has been included. Recent studies have served to

enhance the magnitude scaling factors originally derived by Seed and Idriss (1982). The

results of these studies are presented in Figure 4-4 and indicate that the current scaling factors

are overconservative at earthquake magnitudes less than roughly 7.0 (Arango, 1996).

d)  This plot is appropriate for horizontal sites only. Approximate corrections can be made to the

cyclic stress ratio required to cause liquefaction (CSR l) for sloping sites (Seed and Harder,

1990).

Figure 4-4: Comparison of Earthquake Magnitude Scaling Factors from

Various Sources (Arango, 1996)

e)  The methods for evaluating the liquefaction resistance of soils have been developed for free-

field conditions at horizontal sites where there are no static shear stresses on horizontal

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 planes. This technique is applicable for sites of new construction, yet liquefaction hazard 

studies may also be warranted for existing facilities. Foundation loads imposed by structures

(particularly around the perimeter of the foundation) induce static shear stresses on horizontal

 planes in the soil. These additional stresses will affect the liquefaction susceptibility of the

soil, increasing the hazard in loose soils (relative density ≤ 40%) and potentially decreasing

the liquefaction susceptibility in medium dense and dense soils (relative density ≥ 50%). Theinfluence of foundation stresses on the liquefaction behavior of soils has been investigated by

Rollins and Seed (1990).

The simplified N-based method of liquefaction hazard evaluation has been shown to

 provide reasonable estimates for the occurrence of liquefaction during numerous recent

earthquakes. This agreement with the field case histories has led to its widespread use in

engineering practice.

If the Seed and Idriss method clearly demonstrates the absence of liquefaction hazard,

then this investigation may be sufficient. However, if the occurrence of liquefaction is predicted,

additional seismic hazards should be addressed. These include:

  Hazards associated with lenses of liquefiable soil or by potentially liquefiable layers which

underlie resistant, nonliquefiable capping layers. In situations where few, thin lenses of 

liquefiable soil are identified, the interlayering of liquefiable and resistant soils may serve to

minimize structural damage to light, ductile structures. It may be determined that “life safety”

and/or “serviceability” requirements may be met despite the existence of potentially

liquefiable layers. Ishihara (1985) developed an empirical relation which provides

approximate boundaries for liquefaction-induced surface damage for soil profiles consisting

of a liquefiable layer overlain by a resistant, or protective, surface layer, Figure 4-5. This

relation has been validated by Youd and Garris (1995) for earthquakes with magnitudes

 between 5.3 and 8. In light of the heterogeneous nature of most soil deposits and theuncertainties inherent in the estimation of ground motion parameters, it is recommended that

this method of evaluation be considered for noncritical structures only.

  The potential for lateral ground movements and the effects on foundations and buried 

structures (Bartlett and Youd, 1995).

  The effects of changes in lateral earth pressures on retaining structures (Ebeling and 

Morrison, 1993; Power et al., 1986).

  Estimated total and differential settlements at the ground surface due to liquefaction and 

subsequent densification of the soils (Ishihara and Yoshimine, 1992; Tokimatsu and Seed,1987).

Several of these liquefaction hazards are addressed in the following section.

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Figure 4-5: (A) Relationship between thickness of liquefiable layer and

thickness of overlying layer at sites for which surface manifestation of 

liquefaction has been observed, and (B) guides to evaluation of respective

layer thicknesses (after Ishihara, 1985)

Post-Liquefaction Behavior of Sandy Soils

Post-Liquefaction Volume Change of Sandy Soils The densification of partially-saturated or saturated loose sandy soils due to cyclic loading can result in damaging differential settlement.

This phenomena was graphically demonstrated at Port Island (Port of Kobe) after the 1995

Hyogoken Nanbu earthquake where settlements over much of the island averaged 50 cm, with

maximum settlements of over 1 m in many places. Several methods have been developed for 

estimating the magnitude of earthquake-induced settlements in sandy soils (e.g., Ishihara and 

Yoshimine, 1992; Tokimatsu and Seed, 1987). A simple chart for estimating the volumetric

strain in sandy soils during earthquakes is shown in Figure 4-6..

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Figure 4-6: Post Volumetric Shear Strain for Clean Sands

(Ishihara and Yoshimine, 1992)

Post-Liquefaction Shear Strength In order to evaluate the stability of slopes and embankments,

waterfront retaining structures, and other structures underlain by liquefied soils, the strength of 

the liquefied material must be estimated. Although the condition of initial liquefaction is often

defined as the state at which the effective stress (and therefore the shear strength) is equal to zero,

the soil will mobilize a residual shear strength if it undergoes large shear strains. In two recent

studies, the undrained shear strength of the liquefied sand has been back-calculated from a

number of documented slope failures (Seed and Harder, 1990; Stark and Mesri, 1992). These

reports describe methods for estimating the residual undrained shear strength of sands based on

the SPT penetration resistance of the soil prior to the earthquake, Figure 4-7. The undrained 

strength values obtained from these relations are used in standard total stress stability analysesequivalent to those commonly performed when evaluating the short term stability of 

embankments or foundations on saturated clayey soils.

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Figure 4-7. Relationship between Undrained Critical Strength Ratio and

Equivalent Clean Sand Blow Count (Stark and Mesri, 1992)

Liquefaction-Induced Ground Failures The use of the Seed and Idriss method of evaluating

liquefaction hazard constitutes what may be termed a “triggering” analysis. This term is used to

indicate that the method identifies the CSR required for the surface manifestation of liquefaction.

Soils beneath slopes of as little as 0.2% may experience flow failures subsequent to the onset of 

initial liquefaction. In order to evaluate the seismic performance of pile supported structures,

 breakwaters, pipelines and other structures near slopes, it is necessary to estimate the lateral

deformation that is likely to occur during the design level earthquakes. Both empirical and 

numerical techniques have been developed for this analysis. Youd and his co-workers have

developed simplified procedures for estimating the magnitude of lateral displacements based on

data from numerous case histories (Youd and Perkins, 1987; Bartlett and Youd, 1995). A

regression analysis of field data has resulted in the development of an equation for predicting the

lateral deformation for free-field sites (i.e., in the absence of piles and structures) on gentle

slopes or adjacent to free-faces such as stream banks or dredged channels.

The empirical techniques have been augmented by the results of several numerical studies.

The numerical studies are largely based on the “sliding block” technique (described in Section

4.6) wherein coherent blocks of soil are modeled as moving over a liquefied layer. In this case

the undrained residual (or “steady state”) strength will control the behavior of the soil mass. The

seismic stability of slopes underlain by potentially liquefiable soils can be assessed using

estimated values for the undrained shear strength of the sandy soils.

The transition from the static shear strength to the undrained residual strength requires

several cycles of loading, and it would be advantageous to account for this progressive strength

loss in ground response analyses. Also, the flow behavior of the soil and subsequent

reconsolidation should, theoretically, be modeled in analyses involving soil liquefaction. In light

of this stress-strain-strength behavior, the phenomena of liquefaction-induced ground failure is

complex, and the numerical models used for the evaluation of this hazard have relied on

simplifying assumptions. The sliding block method has been used in conjunction with residual

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undrained strengths for sands to predict the seismic performance of slopes and earth structures

(Baziar et al., 1992; Byrne et al., 1994).

Code Provisions and Factors Of Safety Against Liquefaction

In general building codes do not give extensive guidance for liquefaction apart for theneed for investigating a site for geologic hazards. The AASHTO Standard Specification For 

Highway Bridges (1992) suggests the factor of safety of 1.5 is desirable to establish a reasonable

measure of safety against liquefaction in cases of important bridge sites. While not specifically

stated it is presumed that this is to be used in conjunction with their acceleration maps which

give a 10 percent probability of exceedance in 50 years. Similar recommendations have been

 provided in the CDMG Guidelines for Evaluation and Mitigating Seismic Hazards in California

(1997). The geotechnical panel assembled for the preparation of the CDMG report recommended 

that “If the screening investigation does not conclusively eliminate the possibility of liquefaction

hazards at a proposed project site (a factor of safety of 1.5 or greater), then more extensive

studies are necessary.”

Techniques for Mitigating Liquefaction Hazards

If hazard evaluations indicate that there are potentially liquefiable soils of such extent and 

location that an unacceptable level of risk is presented, then soil improvement methods should be

considered for the mitigation of these hazards. Mitigation must provide suitable levels of 

 protection with regard to the three general types of liquefaction hazard previously noted: (a)

 potential global translational site instability; (b) more localized problems; and (c) failure of 

retaining structures.

Liquefaction remediation must address the specifics of the problem on a case by case

 basis. These specifics include the local site conditions, the type of structure, and the potential for 

flows and settlements. When liquefaction occurs, there can be a potential for extensive lateral

flow slides which can affect a large area, a global site instability. Also there can be local soil

settlements and bearing failures which affect a structure on a local level. Specific types of 

structures can have specific associated problems. Buried structures can become buoyant.

Retaining structures where the backfill has liquefied can experience increased lateral loading and 

deformation.

Potentially suitable methods of mitigation may include the following: removal and 

replacement, dewatering, in-situ soil improvement, containment or encapsulation structures,modification of site geometry, deep foundations, structural systems and, if possible, alternate site

selection. In general, soil improvement methods reduce the liquefaction susceptibility of sandy

soils by increasing the relative density, providing conduits for the dissipation of excess pore

 pressures generated during earthquakes, and/or providing a cohesive strength to the soil. The

effectiveness and economy of any method, or combination of methods, will depend on geologic

and hydrologic factors (e.g.; soil stratigraphy, degree of saturation, location of ground water 

table, depth of improvement, volume of soil to be improved, etc.) as well as site factors (e.g.;

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accessibility, proximity to existing structures, etc.). An overview of the available liquefaction

remediation measures is provided in Table 4-1.

Significant advances have been made in the field of soil improvement over the past

several decades. These developments have been made on two fronts: (a) an increased 

understanding of geotechnical hazards; and (b) the development of innovative constructiontechniques by specialty contractors. Practical information which addresses recent advances in the

mitigation of liquefaction hazards to foundations can be found in numerous recent publications

(Borden et al., 1992; Hryciw, 1995; Kramer and Siddharthan, 1995). Several techniques of 

improvement for liquefiable soils referenced in the geotechnical literature include: (a)

densification (e.g.; vibro-methods, dynamic compaction, deep blast

Table 4-1. Liquefaction Remediation Measures (after Ferritto, 1997B)

Method Principle

Most Suitable

Soil Conditions

or Types

Maximum

Effective

Treatment Depth

Relative

Costs

1)  Vibratory Probe

a)  Terraprobe

 b)  Vibrorods

c)  Vibrowing

Densification by vibration;

liquefaction-induced settlement

and settlement in dry soil under 

overburden to produce a higher 

density.

Saturated or dry

clean sand; sand.

20 m routinely

(ineffective above 3-

4 m depth); > 30 m

sometimes;

vibrowing, 40 m.

Moderate

2)  Vibrocompaction

a)  Vibrofloat

 b)  Vibro-Composer 

system.

Densification by vibration and 

compaction of backfill material

of sand or gravel.

Cohesionless

soils with less

than 20% fines.

> 20 m Low to

moderate

3)  Compaction Piles Densification by displacement of 

 pile volume and by vibration

during driving, increase in

lateral effective earth pressure.

Loose sandy soil;

 partly saturated 

clayey soil; loess.

> 20 m Moderate

to high

4) 

Heavy tamping

(dynamic

compaction)

Repeated application of high-

intensity impacts at surface.

Cohesionless

soils best, other 

types can also be

improved.

30 m (possibly

deeper)

Low

5)  Displacement

(compaction grout)

Highly viscous grout acts as

radial hydraulic jack when

 pumped in under high pressure.

All soils. Unlimited Low to

moderate

6)  Surcharge/buttress The weight of a

surcharge/buttress increases the

liquefaction resistance by

increasing the effective

confining pressures in the

foundation.

Can be placed on

any soil surface.

Dependent on size

of 

surcharge/buttress

Moderate

if vertical

drains are

used 

7)  Drains

a)  Gravel b)  Sand 

c)  Wick 

d)  Wells (for 

 permanent

dewatering)

Relief of excess pore water 

 pressure to prevent liquefaction.(Wick drains have comparable

 permeability to sand drains).

Primarily gravel drains;

sand/wick may supplement

gravel drain or relieve existing

excess pore water pressure.

Permanent dewatering with

 pumps.

Sand, silt, clay. Gravel and sand >

30 m; depth limited  by vibratory

equipment; wick, >

45 m

Moderate

to high

8)  Particulate

grouting

Penetration grouting-fill soil

 pores with soil, cement, and/or 

clay.

Medium to coarse

sand and gravel.

Unlimited Lowest of  

grout

methods

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Method Principle

Most Suitable

Soil Conditions

or Types

Maximum

Effective

Treatment Depth

Relative

Costs

9)  Chemical grouting Solutions of two or more

chemicals react in soil pores to

form a gel or a solid precipitate.

Medium silts and 

coarser.

Unlimited High

10)  Pressure injected 

lime

Penetration grouting-fill soil

 pores with lime

Medium to coarse

sand and gravel.

Unlimited Low

11)  Electrokinetic

injection

Stabilizing chemical moved into

and fills soil pores by electro-

osmosis or colloids in to pores

 by electrphoresis.

Saturated sands,

silts, silty clays.

Unknown Expensive

12)  Jet grouting High-speed jets at depth

excavate, inject, and mix a

stabilizer with soil to form

columns or panels.

Sands, silts, clays. Unknown High

13)  Mix-in-place piles

and walls

Lime, cement or asphalt

introduced through rotating

auger or special in-place mixer.

Sand, silts, clays,

all soft or loose

inorganic soils.

> 20 m (60 m

obtained in Japan)

High

14)  Vibro-replacement

stone and sand 

columns

a)  Grouted 

 b)   Not grouted 

Hole jetted into fine-grained soil

and backfilled with densely

compacted gravel or sand hole

formed in cohesionless soils by

vibro techniques and compaction

of backfilled gravel or sand. For 

grouted columns, voids filled 

with a grout.

Sands, silts, clays. > 30 m (limited by

vibratory

equipment)

Moderate

15)  Root piles, soil

nailing

Small-diameter inclusions used 

to carry tension, shear,

compression.

All soils. Unknown Moderate

to high

16)  Blasting Shock waves an vibrations cause

limited liquefaction,

displacement, remolding, and 

settlement to higher density.

Saturated, clean

sand; partly

saturated sands

and silts after 

flooding.

> 40 m Low

densification, compaction grouting, and compaction piles); (b) drainage to allow rapid 

dissipation of excess pore pressures (e.g.; vibro-replacement and stone columns); and (c)

chemical modification/cementation (e.g.; permeation grouting, jet grouting, and deep mixing).

Additional possible methods of increasing the liquefaction resistance of soils include permanent

dewatering, and removal of loose soils and replacement at a suitable compactive effort.

The most common methods for remediation of liquefaction hazards at open, undeveloped 

sites include (in order of increasing cost): (a) deep dynamic compaction; (b) vibro-compaction;

(c) vibro-replacement, excavation and replacement; and (c) grouting methods. Each of these

techniques results in a significant displacement of soil. On projects involving foundationremediation adjacent to existing structures or buried utilities, ground movements must be

minimized to avoid architectural and structural damage. Several projects with these constraints

have utilized grouting techniques to stabilize potentially liquefiable soils.

Several methods of densification have been used including vibroprobe, vibro-compaction,

dynamic compaction, compaction grouting, and compaction piles. Substitution or replacement of 

soil to improve drainage has been used including vibro-replacement and stone columns.

Techniques like stone columns achieve their effectiveness by replacing liquefiable cohesionless

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soils with stiffer columns of gravel and rock which improves strength and promotes drainage.

Cement grouting, jet grouting and deep mixing have been used as chemical means of 

eliminating/reducing liquefaction potential. Surcharging a site increases liquefaction resistance

 by increasing the effective confining pressures. Table 4  presents a summary of methods used for 

remediation and their relative cost as reported by Professor Whitman (NRC 1985). Navy

facilities on Treasure Island during the 1989 Loma Prieta earthquake can attest to theeffectiveness of remediation. Areas where remediation was done performed well while other 

areas suffered settlements of 6 to 8 inches and lateral spreads. Observation of damage during the

1995 Hyogoken Nanbu (Kobe) Earthquake again confirmed the performance of improved sites.

Preloading, sand drains, sand compaction piles, and vibro-compaction were shown to be

effective.

Method Vertical Settlement

Range

(cm)

Average

(cm)Untreated 25 to 95 42

Preloading 15 to 60 30

Sand drains 0 to 40 15

Sand drains & preloading 0 to 25 12

Vibro-compaction 0 to 5 near 0

Sand compaction piles 0 to 5 near 0

Generally costs increase from dynamic compaction to vibro-compaction to replacement The

measure of effectiveness of a remediation undertaking is the increase in minimum soil density

and specifications usually measure this by the improvement in penetration resistance or laboratory testing. Engineering practice tends to be conservative and factors of safety from 1.5 to

2.0 against liquefaction are often specified. These values may be harder to achieve at the

waterfront in regions of high seismicity.

The contract specifications for ground improvement must include a program for the

verification of soil improvement. Ground improvement specifications typically require that a

minimum soil density (as related to a penetration resistance) be achieved. A level of risk can be

assessed by the factor of safety against liquefaction -- which is defined as the ratio of the CSR 

required for liquefaction of the improved soil to the CSR induced by the design earthquake.

Factors of safety between 1.5 and 2.0 are commonly specified for non-critical structures. It

should be noted that the penetration resistance of recently deposited, or modified, soils increaseswith time after treatment. This phenomena, termed “aging” must be accounted for when

specifying the time between soil improvement and the in-situ verification testing. There is also

evidence that the SPT and CPT tests are not sufficiently sensitive to detect the minor changes in

soil fabric that can significantly increase the liquefaction resistance of the soil. This has led to the

incorporation of lab testing programs in addition to field tests for verification of soil

improvement for liquefaction hazards.

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There are currently very few references on the volume of soil that should be improved to

mitigate liquefaction hazards to structures (JPHRI, 1997; Dickenson and McCullough, 1998). At

horizontal sites, general recommendations for buildings and pile-supported structures may call

for soil improvement to the base of the liquefiable deposit (or to a maximum depth of 15 to 18

meters) and over a lateral extent equal to the thickness of the layer, plus an additional increment

 based on judgment. The uncertainty associated with general recommendations like these iscompounded for sloping sites and areas adjacent to waterfront retaining structures. In these

situations, the volume of soil that may be involved in a lateral spread is difficult to ascertain,

therefore existing recommendations tend to be very conservative. Based on shake table tests, Iai

(1992) has proposed tentative guidelines for the extent of soil improvement required to mitigate

liquefaction hazards behind caissons. These recommendations have been incorporated into the

guidelines for soil improvement adjacent to waterfront retaining structures developed by the

Japan Port and Harbor Research Institute (JPHRI, 1997). Several examples for soil improvement

adjacent to waterfront retaining structures are contained in Figure 4-8.

Very few case studies exist for the seismic performance of improved soil sites. The

effectiveness of the soil improvement at limiting deformations adjacent to foundations and retaining structures will be a function of the strength and duration of shaking. In only a limited 

number of cases have the improved sites been subjected to design-level ground motions. The

cases that have been documented demonstrate a substantial reduction in liquefaction-induced 

ground failures and ground deformations due to the ground treatment (e.g., Iai et al., 1994;

Ohsaki, 1970; Mitchell et al., 1995; Yasuda et al., 1996). However, as noted in several of these

 papers, the ground deformations were not reduced to imperceptible levels. This observation is

especially germane when establishing allowable deformation limits for waterfront retaining

structures, where adjacent gantry cranes and other sensitive components may be damaged by

lateral movements of the walls. In several cases (e.g., flexible retaining structures such as

anchored sheet pile bulkheads and cellular sheet pile bulkheads), these structures may deform

when subjected to strong ground motions despite the utilization of ground treatment. In casessuch as this, the ground treatment would serve to preclude catastrophic failure of the retaining

structures.

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Figure 4-8. Examples of Soil Improvement for Waterfront Retaining

Structures (PHRI, 1997)

One particularly relevant case study involving the waterfront retaining structures at the

Kushiro Port in Japan has been documented by Iai and his co-workers at the Port and Harbour Research Institute (Iai et al., 1994). Kushiro Port is located on the east coast of the island of 

Hokkaido, a region that is prone to large subduction zone earthquakes. Prior to 1993, the port had 

 been subjected to at least two damaging earthquakes (the MJMA  8.1 1952 Tokachi-Oki

Earthquake and the MJMA  7.4 1973 Nemuro-Hanto-Oki Earthquake) and these experiences

appear to have influenced the seismic design criteria subsequently adopted at the port. In the

older portions of the port, the sandy backfill soil adjacent to retaining structures had been left in

its original loose state while, the in the newer sections of the port (constructed as late as 1992), a

soil improvement program was implemented to reduce the liquefaction susceptibility of the fills

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near retaining walls. The remediation project called for the use of various types and degrees of 

soil improvement in the waterfront. In 1993, the port was subjected to strong ground motions

generated during the MJMA 7.8 Kushiro-Oki Earthquake. It is significant to note that while the

waterfront areas at which soil improvement was implemented performed successfully, thereby

allowing the port to continue operation immediately after the earthquake, the waterfront retaining

structures founded in loose soils experienced dramatic liquefaction-induced failures.

Slope Instability Hazards

Introduction

Sloping ground conditions exist throughout ports as natural and engineered embankments

(e.g., river levees, sand or rock dikes, etc., and dredged channel slopes). On-shore and submarine

slopes at ports have been found to be vulnerable to earthquake induced deformations. High water 

levels and weak foundation soils common at most ports can result in slopes which have marginal

static stability and which are very susceptible to earthquake induced failures. In addition towaterfront slopes, several recent cases involving failures of steep, natural slopes along marine

terraces located in backland areas have resulted in damage to coastal ports. Large scale

deformations of these slopes can impede shipping and damage adjacent foundations and buried 

structures thereby limiting port operations following earthquakes.

The most commonly used methods for analysis of slope stability under both static and 

dynamic conditions are based on standard rigid body mechanics and limit equilibrium concepts

that are familiar to most engineers. The development and application of these techniques are

introduced in most geotechnical engineering textbooks and in numerous design manuals.

Therefore, they will not be described here. Instead, this section will introduce the strengths and 

limitations of these analysis techniques as applied at port facilities.

Pseudostatic Methods of Analysis

Standard, limit equilibrium methods for analyzing the static stability of slopes are

routinely used in engineering practice. The use of these design tools have several advantages in

 practice: (a) the techniques are familiar to most engineers; (b) requisite input includes standard 

geotechnical parameters that are obtained during routine foundation investigations; and (c) the

methods have been coded in very straightforward, efficient computer programs that allow for 

sensitivity studies to be made of various design options.

For use in determining the seismic stability of slopes, limit equilibrium analyses are

modified slightly with the addition of a permanent lateral body force which is the product of a

seismic coefficient  and the mass of the soil bounded by the potential slip. The seismic coefficient

(usually designated as k h, Nh) is specified as a fraction of the peak horizontal acceleration, due to

the fact that the lateral inertial force is applied for only a short time interval during transient

earthquake loading. Seismic coefficients are commonly specified as roughly1/3 to

1/2 of the peak 

horizontal acceleration value (Seed, 1979; Marcuson, et al., 1992).

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In most cases involving soils which do not exhibit considerable strength loss after the

 peak strength has been mobilized, common pseudostatic rigid body methods of evaluation will

generally suffice for evaluating the stability of slopes. These methods of evaluation are well

established in the technical literature (Kramer, 1996). Although these methods are useful for 

indicating an approximate level of seismic stability in terms of a factor of safety against failure,

they suffer from several potentially important limitations. The primary disadvantages of  pseudostatic methods include: (a) they do not indicate the range of slope deformations that may

 be associated with various factors of safety; (b) the influence of excess pore pressure generation

on the strength of the soils is incorporated in only a very simplified, “decoupled” manner; (c)

 progressive deformations that may result due to cyclic loading at stresses less than those required 

to reduce specific factors of safety to unity are not modeled; (d) strain softening behavior for 

liquefiable soils or sensitive clays is not directly accounted for; and (e) important aspects of soil-

structure interaction are not evaluated.

Limited Deformation Analysis

In most applications involving waterfront slopes and embankments, it is necessary to

estimate the permanent slope deformations that may occur in response to the cyclic loading.

Allowable deformation limits for slopes will reflect the sensitivity of adjacent structures,

foundations and other facilities to these soil movements. Enhancements to traditional

 pseudostatic limit equilibrium methods of embankment analysis have been developed to estimate

embankment deformations for soils which do not lose appreciable strength during earthquake

shaking (Ambraseys and Menu, 1988; Makdisi and Seed, 1978; Jibson, 1993).

Rigid body, “sliding block” analyses, which assume the that soil behaves as a rigid,

 perfectly plastic material, can be used to estimate limited earthquake-induced deformations. The

technique, developed by Newmark (1965) and schematically illustrated in Figure 4-9, is based onsimple limit equilibrium stability analysis for determining the critical, or yield, acceleration

which is required to bring the factor of safety against sliding for a specified block of soil to unity.

The second step involves the introduction of an acceleration time history. When the ground 

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a) b)

Figure 4-9. Elements of Sliding Block Analysis, A) Hynes-Griffin and

Franklin, 1984, B) after Wilson and Keefer, 1985

motion acceleration exceeds the critical acceleration (acrit, ay) the block begins to move down

slope. By double integrating the area of the acceleration time history that exceeds acrit, the relative

displacement of the block is determined. A simple spreadsheet routine can be used to perform

this calculation (Jibson, 1993). Numerical studies based on this method of analysis have lead to

the development of useful relationships between ground motion intensity and the seismically-

induced deformations (e.g. Ambraseys and Menu, 1988; Makdisi and Seed, 1978; Jibson, 1993).

The relationship proposed by Makdisi and Seed is shown in Figure 4-10.

The amount of permanent displacement depends on the maximum magnitude and duration of the earthquake. The ratio of maximum acceleration to yield acceleration of 2.0 will

result in block displacements of the order of a few inches for a magnitude 6 1/2 earthquake and 

several feet for a magnitude 8 earthquake. It should be noted that significant pore pressure

increases may be induced by earthquake loading in saturated silts and sands. For these soils a

 potential exists for a significant strength loss. For dense saturated sand, significant undrained 

shear strength can still be mobilized even when residual pore pressure is high. For loose sands,

the residual undrained strength which can be mobilized after high pore pressure build-up is very

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low and is often less than the static undrained shear strength. This may result in flow slides or 

large ground deformations.

Given that the sliding block analyses are based on limit equilibrium techniques, they

suffer from many of the same deficiencies previously noted for pseudostatic analyses. One of the primary limitations with respect to their application for submarine slopes in weak 

Figure 4-10. Empirical Relationships between Permanent Displacement of 

sliding Block and ratios of accelerations (after Makdisi and Seed, 1978)

soils is that strain softening behavior is not directly accounted incorporated in the analysis. The

sliding block methods have, however, been applied for liquefiable soils by using the post-

liquefaction undrained strengths for sandy soils. A recent method proposed by Byrne et al. (1994)

has been developed for contractive, collapsible soils which are prone to liquefaction.

Advanced Numerical Modeling of Slopes

In situations where the movement of a slop impacts adjacent structures, such as pile

supported structures embedded in dikes, buried lifelines and other soil-structure interaction

 problems, it is becoming more common to rely on numerical modeling methods to estimate the

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range of slope deformations which may be induced by design level ground motions (Finn, 1990).

The numerical models used for soil-structure interaction problems can be broadly classified 

 based on the techniques that are used to account for the deformations of both the soil and the

affected structural element. In many cases the movement of the soil is first computed, then the

response of the structure to these deformations is determined. This type of analysis is termed 

uncoupled , in that the computed soil deformations are not affected by the existence of embedded structural components. A common enhancement to this type of uncoupled analysis includes the

introduction of an iterative solution scheme which modifies the soil deformations based on the

response of the structure so that compatible strains are computed. An example of this type of 

analysis is drag loading on piles due to lateral spreading. In an uncoupled analysis the ground 

deformations would be estimated using either an empirical relationship (e.g. Bartlett and Youd 

1995) or a sliding block type evaluation (e.g. Byrne et al., 1994) as discussed in Section 0 and 0.

Once the pattern of ground deformations has been established a model such as LPILE (Reese and 

Wang, 1994) can be used to determine the loads in the deformed piles. In addition, modifications

can be made to the p-y curves to account for the reduced stiffness of the liquefied soil (O’Rourke

and Meyerson, 1994). The lateral spread displacement is forced onto the p-y spring (i.e., drag

loading).

In a coupled   type of numerical analysis the deformations of the soil and structural

elements are solved concurrently. Two-dimensional numerical models such as FLUSH (Lysmer 

et al., 1975), FLAC (Itasca, 1995), DYSAC2 (Muraleetharan et al., 1988), and LINOS (Bardet,

1990) have been used to model the seismic performance of waterfront components at ports (e.g.,

Finn, 1990; Roth et al., 1992; Werner, 1986; Wittkop, 1994). The primary differences in these

numerical analyses include; (a) the numerical formulation employed (e.g., FEM, FDM, BEM),

(b) the constitutive model used for the soils, and (c) the ability to model large, permanent

deformations. Each of the methods listed have been useful in evaluating various aspects of 

dynamic soil-structure interaction.

Advanced numerical modeling techniques are recommended for soil-structure interaction

applications, such as estimating permanent displacements of slopes and embankments with pile

supported wharves. The primary advantages of these models include: (a) complex embankment

geometries can be evaluated, (b) sensitivity studies can be made to determine the influence of 

various parameters on the seismic stability of the structure, (c) dynamic soil behavior is much

more realistically reproduced, (d) coupled analyses which allow for such factors as excess pore

 pressure generation in contractive soils during ground shaking and the associated reduction of 

soil stiffness and strength can be used, (e) soil-structure interaction and permanent deformations

can be evaluated. Disadvantages of the numerical analyses methods include: (a) the engineering

time required to construct the numerical model can be extensive, (b) numerous soil parameters

are often required, thereby increasing laboratory testing costs (the number of soil propertiesrequired is a function of the constitutive soil model employed in the model), (c) very few of the

available models have been validated with well documented case studies of the seismic

 performance of actual retaining structures, therefore the level of uncertainty in the analysis is

often unknown.

Mitigation of Seismic Hazards Associated with Slope Stability

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Remedial strategies for improving the stability of slopes have been well developed for 

 both onshore and submarine slopes. Common techniques for stabilizing slopes include: (a)

modifying the geometry of the slope; (b) utilization of berms; (c) soil replacement (key trenches

with engineered fill); (d) soil improvement; and (e) structural techniques such as the installation

of piles adjacent to the toe of the slope. Constraints imposed by existing structures and facilities,

and shipping access will often dictate which of the methods, or combinations of methods, areused.

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 References

Ambraseys, N.N., and Menu, J.M. (1988). Earthquake-Induced Ground Displacements,

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Arango, I. (1996). Magnitude Scaling Factors for Soil Liquefaction Evaluations, Journal of 

Geotechnical Engineering, ASCE, Vol. 122, No. 10, pp. 929-936.

Bardet, J.P. (1990). LINOS - A Non-Linear Finite Element Program for Geomechanics and 

Geotechnical Engineering, Civil Engineering Dept., Univ. of Southern California.

Bartlett, S.F., and Youd, T.L. (1995), Empirical Prediction of Liquefaction-Induced Lateral

Spread, Journal of the Geotechnical Engineering, ASCE, Vol. 121, No. 4, pp. 316-329.

Baziar, M.H., Dobry, R., and Elgamal, A-W.M. (1992), Engineering Evaluation of Permanent

Ground Deformations Due to Seismically-Induced Liquefaction, Technical Report NCEER-92-0007, National Center for Earthquake Engineering Research, SUNY at Buffalo, NY.

Borden, R.H., Holtz, R.D., and Juran, I., editors (1992), Grouting, Soil Improvement and 

Geosynthetics, ASCE Geotechnical Special Publication No. 30, , ASCE, NY,NY, 1453 p.

Byrne, P.M., Jitno, H., Anderson, D.L., and Haile, J. (1994). A Procedure for Predicting Seismic

Displacements of Earth Dams, Proc. of the 13th International Conference on Soil Mechanics and 

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Dickenson, S.E., and McCullough, N.J. (1998) Mitigation of Liquefaction Hazards to Anchored 

Sheet Pile Bulkheads. To be published in the Proceeding of ASCE Ports ’98 . ASCE, March

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Egan, J.A., Hayden, R.F., Scheibel, L.L., Otus, M., and Serventi, G.M. (1992). Seismic Repair at

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Finn, W.D.L. (1990). Analysis of Deformations, Stability and Pore Water Pressures in Port

Structures, Proc. of the POLA Seismic Workshop on Seismic Engineering, Port of Los Angeles,

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Geomatrix (1995). Seismic Design Mapping for the State of Oregon, Report for the Oregon

Department of Transportation, Salem, Oregon, Personal Services Contract 11688.

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Tanaka, Y., and Kato, M. (1994). Effects of Remedial Measures Against Liquefaction at 1993

Kushiro-Oki Earthquake. Proceedings from the Fifth U.S.-Japan Workshop on Earthquake

 Resistant Design of Lifeline Facilities and Countermeasures Against Soil Liquefaction.

Technical Report NCEER-97-0026. O’Rourke and Hamada (eds.). National Center for 

Earthquake Engineering Research, State University of New York at Buffalo.754 p.

Inagaki, H., Iai, S., Sugano, T., Yamazaki, H., and Inatomi, T. (1996). Performance of Caisson

Type Quay Walls at Kobe Port, Soils and Foundations, Special Issue on Geotechnical Aspects of 

the January 17, 1995 Hyogoken Nanbu Earthquake, January, pp. 119-136.

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